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Teall, Oliver, Pilegis, Martins, Davies, Robert, Sweeney, John, Jefferson, Tony, Lark, Robert and
Gardner, Diane 2018. A shape memory polymer concrete crack closure system activated by
electrical current. Smart Materials and Structures 27 (7) , 075016. 10.1088/1361-665X/aac28a file
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A shape memory polymer concrete crack closuresystem activated by electrical currentTo cite this article: Oliver Teall et al 2018 Smart Mater. Struct. 27 075016
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A shape memory polymer concrete crack
closure system activated by electrical
current
Oliver Teall1,4
, Martins Pilegis2, Robert Davies
2, John Sweeney
3,
Tony Jefferson2, Robert Lark
2and Diane Gardner
2
1Costain Group PLC., Costain House, Vanwall Business Park, Maidenhead, Berkshire SL6 4UB, United
Kingdom2Cardiff University School of Engineering, Queen’s Buildings, 14-17 The Parade, Cardiff CF24 3AA,
United Kingdom3Bradford University School of Engineering, Bradford, West Yorkshire BD7 1DP, United Kingdom
E-mail: [email protected]
Received 10 February 2018, revised 1 May 2018
Accepted for publication 4 May 2018
Published 31 May 2018
Abstract
The presence of cracks has a negative impact on the durability of concrete by providing paths for
corrosive materials to the embedded steel reinforcement. Cracks in concrete can be closed using
shape memory polymers (SMP) which produce a compressive stress across the crack faces. This
stress has been previously found to enhance the load recovery associated with autogenous self-
healing. This paper details the experiments undertaken to incorporate SMP tendons containing
polyethylene terephthalate (PET) filaments into reinforced and unreinforced 500×100×100mm
structural concrete beam samples. These tendons are activated via an electrical supply using a nickel-
chrome resistance wire heating system. The set-up, methodology and results of restrained shrinkage
stress and crack closure experiments are explained. Crack closure of up to 85% in unreinforced
beams and 26%–39% in reinforced beams is measured using crack-mouth opening displacement,
microscope and digital image correlation equipment. Conclusions are made as to the effectiveness of
the system and its potential for application within industry.
Keywords: shape memory polymer, self-healing concrete, durability, concrete
(Some figures may appear in colour only in the online journal)
1. Introduction
Cracking in reinforced concrete structures is a common
phenomenon, with a variety of causes. It is generally accepted
that the presence of cracks has a negative impact on the
durability of concrete by providing paths for corrosive
materials to the embedded steel reinforcement (Van Breugel
2007, Van den Heede et al 2014).
Concrete is capable of natural or ‘autogenous’ healing in
certain conditions. The width of the crack is a fundamental
factor which affects the self-healing process (De Rooij
et al 2013). Methods to restrict crack widths within concrete
commonly include the introduction of steel or polymer fibres.
Steel cord fibres have been found to corrode within concrete
cracks, making them less effective as a crack control mech-
anism for self-healing purposes (Homma et al 2009).
Engineered cementitious composite (ECC) materials contain
polymer fibres, typically 2% by volume, which help to reduce the
appearance of single cracks and distribute any micro-cracks
throughout the material, resulting in a smaller average crack
Smart Materials and Structures
Smart Mater. Struct. 27 (2018) 075016 (12pp) https://doi.org/10.1088/1361-665X/aac28a
4Author to whom any correspondence should be addressed.
Original content from this work may be used under the terms
of the Creative Commons Attribution 3.0 licence. Any
further distribution of this work must maintain attribution to the author(s) and
the title of the work, journal citation and DOI.
0964-1726/18/075016+12$33.00 © 2018 IOP Publishing Ltd Printed in the UK1
width compared to concrete reinforced with conventional steel
bars. ECC can be designed to reduce average crack widths to as
low as 0.03mm (Li et al 1998, Yang et al 2009, De Rooij
et al 2013).
However, A typical ECC mix contains 570 kg m−3
portland cement and 684 kg m−3 pulverised fly ash (De Rooij
et al 2013), high proportions compared to conventional
reinforced concrete. While contributing to self-healing by on-
going hydration, this results in a higher level of embodied
CO2 and increased material cost compared to more conven-
tional mixes. Given the UK government’s aim, set out in
‘Construction 2025’ (Government 2013), to achieve a 50%
reduction in greenhouse gas emissions and 33% lower initial
construction costs in the built environment by 2025, a solu-
tion which produces the same benefits as ECC while reducing
the cement content would be desirable.
The application of a compressive stress across the crack
faces has been shown to improve strength recovery with auto-
genous healing (Neville 2012). Heide and Schlangen (2007)
investigated the autogenous healing of submerged concrete
samples subjected to an applied compressive stress following
three-point bending to induce cracks. They found that with a
compressive stress of 1 Nmm−2 on the crack face, the stiffness
and strength of the concrete specimen was recovered, while
without this stress the strength recovery was minor.
Shape memory polymers (SMP) have been proposed as a
method to apply this compressive stress within concrete struc-
tures (Jefferson et al 2010, Isaacs et al 2013). SMP are materials
which respond to external stimuli by altering their physical
shapes. The most common SMP are thermoresponsive, with
direct heat used as the stimulus for shape change (Xie 2011). If
an amorphous SMP, which has previously been oriented by
drawing, is heated above its glass transition temperature (Tg)
while restrained, a restrained shrinkage stress is produced which
can be utilised to close cracks in concrete.
Over the last decade, potential biomedical applications have
contributed to a large increase in SMP interest, examples of
which include active medical devices such as a laser-activated
device for the mechanical removal of blood clots (Xie 2011,
Behl and Lendlein 2007). However, uses for SMPs have
also been found in a wide range of other areas, such as smart
textiles, high performance water vapour permeability materials,
heat shrinkable electronics packaging, sensors and actuators,
self-deployable structures in spacecraft and micro-systems
(solution, emulsion, film, foam and bulk) (Meng and Hu 2009).
Proof of concept experiments carried out by Jefferson et al
(2010) and Isaacs et al (2013) used drawn polyethylene ter-
ephthalate (PET) tape to close cracks in prismatic hollow
mortar beam specimens, making use of the shrinkage stress
generated by the material’s retraction upon heating. Commer-
cially available PET shrink tite samples were placed within
prismatic hollow mortar beam samples, 25×25×150mm in
dimension with a 10×10mm void running through the cen-
tre. The mix contained water:cement:sand contents of
306 kgm−3:510 kgm−3:1530 kgm−3(or 0.6:1:3 by weight).
Layers of strips, with a total cross-sectional area of 20.7mm2,
were placed through the void and externally anchored. A three-
stage test was carried out, consisting of three-point loading to
form a 0.3 mm width crack, dry heating in an oven at 90 °C to
activate the PET specimens and a second round of three-point
loading to failure. Conclusions from these tests were that
activation of the tendons closed the cracks and applied a pre-
stress of between 1.5 and 2MPa on the mortar beam specimens
(Jefferson et al 2010).
However, these experiments were carried out at a small
scale within a laboratory setting, and use mortar beams rather
than structural concrete. As suggested by Isaacs (2011), larger
specimens may require polymers which generate larger
shrinkage stresses for lower concrete volume replacement. A
suitable activation technique for the polymers is also required.
Jefferson et al (2010) suggest that the use of ovens to activate
the PET shrink tite is regarded as a temporary solution by
stating that ‘It is envisaged that in later versions of the system,
activation will be achieved via an electrical supply’.
This paper describes the experimental work undertaken to
incorporate SMP tendons, made from bundled PET filaments,
into 500×100×100mm structural concrete beams. The
beams were cracked in three-point bending and the tendons
activated via an electrical current using a resistance wire heating
system. Crack closure was measured using a crack-mouth
opening displacement (CMOD) gauge and microscope images,
with digital image correlation (DIC) used for qualitative con-
firmation. Testing undertaken to determine the restrained
shrinkage stress generated by the filaments is described in detail
in a previous paper by the author (Teall et al 2017).
The experiments in this paper demonstrate the feasibility
of a SMP concrete crack closure system activated by electrical
current, a novel approach to reducing the width of concrete
cracks which has not previously been described.
2. PET tendons
The PET tendons are shown in figure 1 below. Each tendon
was 400 mm in length and contained 200 PET filaments. The
filaments were manufactured at Bradford University using a
commercial grade of PET (Dow Lighter C93). The manu-
facturing and die-drawing process for the filaments is
described in detail in Teall et al (2017). The final diameter of
the filaments was 0.9 mm, with a draw ratio of 4.0.
To form the tendons, the filaments were wound around
two pegs on a bespoke steel rig, the schematic for which is
shown in figure 2. M10 (10 mm diameter) stainless steel wire
rope clips were used to grip the filaments at either end
(figure 3) and an injection-moulded polylactic acid (PLA)
sleeve placed around the tendon to de-bond and protect the
filaments from the surrounding concrete. A steel plate was
placed between the wire rope clip and the sleeve. The area of
the plate extended beyond the edge of the PLA sleeve to
provide a flat surface to transfer the compressive stress from
the tendon to the concrete, and from the tendon to the testing
rig during the restrained shrinkage tests.
PTFE-coated nickel-chrome (nichrome) wire was wrap-
ped around the filaments to provide the heat for activation
when passing a current through the wire. The PET filaments
were coated in silicone grease to improve heat transfer. A coil
2
Smart Mater. Struct. 27 (2018) 075016 O Teall et al
spacing of 10 mm was used for the nichrome wires. Two
thermocouples were incorporated into each tendon, as shown
in figure 4, to measure temperature during activation. The
first, ‘internal’ thermocouple was placed near the centre of the
filament bundle. The second ‘external’ thermocouple was
placed on the surface of the filaments, between two heating
coils.
3. Experiment set-up
3.1. Restrained shrinkage test
To test the restrained shrinkage stress produced by each
tendon, they were placed within a steel grip arrangement as
shown in figure 5. The steel grips were connected by a
threaded bar and nut to a 5 kN full bridge strain gauge to
measure load. The steel plates at either end of the tendon
rested on the steel grips, providing a surface for transferring
load. Before testing, the tendons were loaded to a pre-stress of
2 MPa by increasing the distance between the two steel grips
Figure 1. Shape memory PET tendon concrete crack closure system (a) schematic (b) photo.
Figure 2. Schematic of bespoke steel rig for making filamenttendons.
Figure 3. (a) Wire rope clips; (b) wire rope clip on PET filaments.
3
Smart Mater. Struct. 27 (2018) 075016 O Teall et al
using the threaded bar. Each end of the nichrome heating wire
was connected to a length of low-resistance copper wire,
which was in turn attached to an Ashley-Edison AE-130-LB
variable transformer connected to a 230 V mains supply,
allowing the input voltage to be manually varied.
3.2. Concrete crack closure tests
The concrete beam samples measured 500×100×100 mm
and were cast using steel moulds. Each of the test beams
contained one PET tendon. The reinforced samples also
contained two 6 mm diameter high tensile ribbed steel bars.
The set-up of the reinforced samples prior to casting is shown
in figure 6 and the cross-sections of unreinforced and rein-
forced beam samples are shown in figure 7.
The concrete mix design shown in table 1 was mixed
using a Belle Premier 200XT mixer and compacted in three
layers on a vibrating table. Coarse aggregate consisted of
4–10 mm diameter limestone. Fine aggregate used was
0–4 mm diameter natural sea-dredged sand from Swansea
Wharf and 0–2 mm diameter limestone dust sieved from
0–4 mm diameter limestone aggregates. All of the limestone
was sourced from Cornelly Drystone Quarry in Porthcawl,
South Wales. The cement used was CEM II/B-V 32, 5R with
21%–35% Fly Ash content, consistent with BS EN 197-1
classifications. This was sourced from Lafarge Tarmac.
The target consistency of the fresh concrete, measured
using the slump test to BS EN 206:2013 (BSI 2013), was S3
(100–150 mm). Adomast Adoflow Extra plasticiser was
added to the mix after all other constituents to reach the target
slump range, in amounts not exceeding 1.75 kg m−3.
After air-curing for 24 h, the specimens were submerged
in water at 20 °C for 5 d. They were then removed from the
water, a 5 mm deep notch was cut as a crack-inducer at mid-
span and knife edges attached to measure CMOD. The beams
were left for a further 24 h to dry in ambient conditions,
before testing at 7 d after casting. The beam sample references
and descriptions are given in table 2 and the full testing
methodology described in table 3.
4. Experimental methodology
4.1. Restrained shrinkage test
The input voltage was raised until 90 °C was recorded on the
internal thermocouple. The temperature was then regulated by
manually switching the system on and off to keep the reading
within 5 °C of this target temperature for 15 min. This caused
the temperature measured by the external thermocouple to
fluctuate between 80 °C and 90 °C due to a higher rate of heat
Figure 4. Position of thermocouples within PET filament tendon.
Figure 5. Shape memory PET tendon testing rig arrangement.
Figure 6. Reinforced test sample set-up prior to casting.
4
Smart Mater. Struct. 27 (2018) 075016 O Teall et al
loss on the external face of the tendon. The electricity supply
was then turned off, allowing the tendon to cool to room
temperature. The load and temperature were measured at a
rate of 2 Hz using the load cell and thermocouples respec-
tively. This test was repeated with three tendons.
4.2. Concrete crack closure tests
Table 3 describes the testing methodology for the PET tendon
in unreinforced and reinforced beam samples. Cube samples
were also tested at 7 and 28 d for compressive strength. The
three-point bending set-up is shown in figure 8.
4.3. Crack width and strain measurements
Images of the cracks were taken on both sides of the beam
samples using a Veho Discovery VMS-001 USB microscope
along with a graduated scale. Images were taken at peak load
and after unloading on both the initial and reload cycles.
Additional images were taken immediately after activation of
the tendon. From each image, three crack width measure-
ments 0.75 mm apart were made perpendicular to the crack
face using ImageJ software (Schneider et al 2012), as illu-
strated in figure 9. Average crack width measurements from
both sides of the beam were used for comparison.
Displacement and strain were also measured using a DIC
system, consisting of two LaVision Imager X-lite 8 M CCD
Figure 7. Cross-section of reinforced and unreinforced test beams.
Table 1. Nominal concrete mix design (saturated surface dry quantities).
Cement
(kg m−3)
w/cratio
Coarse aggregate
(kg m−3)
Fine aggregate
(kg m−3)
Limestone dust
(kg m−3)
Water
(kg m−3)
Admixture
(kg m−3)
400 0.48 990 648 162 192a 1.75b
a
Nominal SSD quantity, actual added water adjusted according to aggregate water content.b
Plasticiser was added to reach the target slump range in amounts not exceeding 1.75 kg m−3.
Table 2. Specimen descriptions.
Sample Series Description
PET A Unreinforced 500×100×100 mm concrete beam containing 1 PET tendon
PET B Unreinforced 500×100×100 mm concrete beam containing 1 PET tendon
PET C Unreinforced 500×100×100 mm concrete beam containing 1 PET tendon
Control (x3) Unreinforced 500×100×100 mm concrete beam
PET A-r Reinforced 500×100×100 mm concrete beam containing 1 PET tendon and 2 no. 6 mm diameter
high tensile steel bars
PET B-r Reinforced 500×100×100 mm concrete beam containing 1 PET tendon and 2 no. 6 mm diameter
high tensile steel bars
PET C-r Reinforced 500×100×100 mm concrete beam containing 1 PET tendon and 2 no. 6 mm diameter
high tensile steel bars
CONT A-r Reinforced 500×100×100 mm concrete beam containing 2 no. 6 mm diameter high tensile steel bars
CONT B-r Reinforced 500×100×100 mm concrete beam containing 2 no. 6 mm diameter high tensile steel bars
CONT C-r Reinforced 500×100×100 mm concrete beam containing 2 no. 6 mm diameter high tensile steel bars
Cubes (x6) Unreinforced and reinforced 100×100×100 mm cube samples for compression tests
5
Smart Mater. Struct. 27 (2018) 075016 O Teall et al
cameras and DaVis analysis software. A speckle pattern (to
create unique subsets) was painted (using spray paint) on to
the side of all beams containing PET tendons and one control
beam from both the unreinforced and reinforced beam series.
A typical example of this pattern is shown in figure 10.
Images were taken every 30 s during loading, activation and
reloading using the DIC system. These images were then
Table 3. Testing methodology for PET tendons in unreinforced and reinforced beam samples.
Day from
casting Action Description
0 Casting of all beams Beams cast and left to cure in air
1 Submersion Beams removed from steel moulds and submerged in water at 20 °C
6 Preparation for testing Removal from water, given 5 mm notch and knife edges
7 Initial three-point bend testing of
all beams
Unreinforced beams (PET A-C and controls)
Loaded to 0.5 mm CMOD at a rate of 0.001 mm s−1(CMOD controlled). At 0.5 mm
the CMOD was held constant while microscope images were taken of both sides of
the beam (at crack location). The beams were then unloaded and additional
microscope images taken
Reinforced beams (PET A-r, B-r, C-r and CONT A-r, B-r, C-r)
Loaded to 15 kN at a rate of 0.001 mm s−1(CMOD controlled). At 15 kN the CMOD
was held constant while microscope images were taken of both sides of the beam
(at crack location). The beams were then unloaded and additional microscope
images taken.
7 Activation of PET tendons Immediately after initial testing, the tendons were activated by passing a current
through the nichrome wire. The voltage was increased until either a temperature
plateau of 90 °C was reached on the internal thermocouple, or the wire temperature
reached 150 °Ca. This voltage was then kept constant for 1 hb before turning off the
electricity supply and allowing the tendon to cool to room temperature.
Cont B-r was also left for an hour as a control to test the effect of the reinforcing steel
on the gradual crack closure over this time period
7 Retesting of beams in three-point
bending
Unreinforced beams (PET A-C)
Reloaded to 1 mm CMOD at a rate of 0.001 mm s−1, additional microscope images
taken, then unloaded
Reinforced beams (PET A-r, B-r, C-r and CONT A-r, B-r, C-r)
Reloaded to 15 kN at a rate of 0.001 mm s−1, additional microscope images taken,
then unloaded. For PET A-r, PET C-r and Cont B-r two additional reloads were
completed to investigate the cyclic loading effect. Cont A-r was also reloaded a
second time before a machine error caused the failure of the beam. Cont C-r was
only reloaded once
a
A separate study on the activation system showed that the temperature on the outside of the sleeve temperature never exceeded 100oC and from this study it is
was concluded that the heating of the tendon would not cause any damage to the surrounding concrete matrix.b
The temperature was not regulated as the thermocouple readings reached a plateau using voltage control alone.
Figure 8. Three-point bend test set-up (dimensions in mm).
Figure 9. Crack measurement using microscope.
6
Smart Mater. Struct. 27 (2018) 075016 O Teall et al
post-processed in the DaVis software to track movement over
time, generating 2D strain images to visualise crack closure.
5. Results and discussion
5.1. Polymer tendon shrinkage stress
Figure 11 shows the graph of stress and temperature against
time from the restrained shrinkage stress test 1 on the PET
tendons. This is typical of all of the tests undertaken. Table 4
shows the peak and final restrained shrinkage stress values for
all three tendons. The final stress is defined as the stress at
30 °C after cooling.
In all three tests, as the temperature was raised above Tg(60 °C–80 °C) the restrained shrinkage stress rapidly
increased, up to a maximum of 19.93–21.39MPa after 15 min
at 90 °C. Upon cooling, there was a gradual drop in shrinkage
stress, attributed to a decrease in the entropic state of the
molecular chains, resulting in a final stress plateau at between
17.76 and 19.19MPa. The coefficient of variation for both the
peak and final stress, shown in table 4, was 3.2% and 3.6%
respectively. The method of manufacture therefore produces
repeatable peak and final stresses in the manufactured
tendons.
5.2. Compressive strength and consistency class
Table 5 shows the compressive strength results from the cube
samples. All of the strength results were above the target
characteristic compressive strength of 45 Nmm−2, with a
mean value exceeding 49 Nmm−2. The mix therefore con-
forms to the strength class C35/45 based on EN 206:2013 in
terms of both strength and variability. The measured slump of
the mix was 150 mm, falling into consistency class S3
(BSI 2013).
5.3. Unreinforced beams
Figure 12 shows typical load-CMOD curves for the unrein-
forced test and control samples and figure 13 the computed
stress distribution diagrams at various stages of the experi-
ment. These were computed assuming that concrete is a no-
tension material, Engineering Beam theory is applicable and
that the tendon is unbonded between anchorages. Comparing
the initial loading peaks (point 1) gives an average reduction
Figure 10. Applied speckle pattern for DIC system.
Figure 11. Restrained shrinkage stress test on PET tendon—test 1.
Table 4. Peak and final restrained shrinkage stress results from PETtendon tests.
Test
Peak
stress (MPa)
Final
stress (MPa)
CoV
peak (%)
CoV
final (%)
1 21.39 19.19
2 19.93 17.83 3.2 3.6
3 20.11 17.76
Table 5. Compressive strength results from cube samples (average of3 cubes).
Age (days) Compressive strength (N mm−2) CoV (%)
7 36.96 2.35
28 51.17 0.55
Figure 12. Load-CMOD curves for unreinforced test and controlbeams—typical graph.
7
Smart Mater. Struct. 27 (2018) 075016 O Teall et al
in peak load of 6.2% for the PET samples (4.78 kN compared
to 5.10 kN for the controls), attributed to the replacement of
concrete with the PET filament tendon, which take up
approximately 1.8% of the cross-sectional area. As illustrated
by figure 13, at point 1 the un-cracked concrete is assumed to
be linear elastic and the polymer tendon to be subject to stress
SP1, which is the stress associated with its elongation between
anchor points at this load level. This results in a stress of
0.5 N mm−2 in the polymer for the typical sample shown in
figure 12.
The shape of the response diagram between points 1 and
2 results from a loss of stiffness due to concrete cracking and
an increase in the tendon force, the latter being caused by the
extension of the tendon between anchor points due to the
overall increase in flexural displacements of the beam.
Assuming that concrete carries no-tension after cracking
initiates, the stress distribution at point 2 may be assumed to
be that shown in figure 13(a). Using the idealised model, the
stress in the polymer at this load point may be calculated as
follows;
=⋅
( )SM
Z A, 1P2
2
P
where M is the applied moment, Z is the lever arm, AP is the
area of the polymer and the numbered subscript denotes the
loading point.
With an applied load of 1.77 kN (for the typical
graph shown), SP2 is 21.6 N mm−2. From previous
experiments detailed within Teall et al (2017), this is known
to be well within the polymer’s elastic zone.
Upon unloading (point 3), a residual CMOD of
0.195 mm to 0.231 mm was recorded. Such residual dis-
placements are typical in concrete fracture tests and are nor-
mally attributed to a mismatch of asperities and local friction
between opposing crack faces(van Mier 1996). The fact that
the residual CMOD of the PET beam is significantly lower
than that of the control sample in figure 12 is due to the load
in the tendon at this point (induced from tendon extension via
the residual crack opening), which exerts an eccentric com-
pressive load to the beam that acts to close the crack. It is also
noted that the loading rig was only capable of applying
downward vertical forces to the beam, which provides zero
the lower bound for P shown in figure 12.
Upon activation (point 3 to point 4), the compressive
stress produced by the tendon caused a reduction in CMOD,
as the crack faces were drawn back together. Figure 13 shows
the idealised stress profile of the beam at this stage, which is
made up of the stresses due to the applied axial load and
moment from the tendon, it being assumed the beam can
accommodate nominal tensile stresses in the upper part of the
beam. Table 6 shows the percentage crack closure of each of
the test samples. This closure was measured both by the
change in CMOD readings and from microscope measure-
ments of the induced crack before and after the activation
process. The recorded CMOD values were larger than the
microscope measurements due to the position of the gauge,
approximately 4 mm below the bottom surface of the beam.
The difference between the two measurements was
0.035–0.063 mm before activation, reducing to 0–0.007 mm
after activation as the crack faces were pulled back together.
The variation in measurements between beams is caused by
differing crack width, variations in actual notch depth and
variation in the exact position of microscope measurements
relative to the CMOD gauge.
The reduction in difference between measurements upon
activation is due to a decrease in beam curvature caused by
the compressive force from the tendon. It is therefore to be
expected that the crack closure measured by change in
CMOD will be slightly higher than those measured by
microscope images.
Activation of the polymer tendon significantly increased
the stiffness of the beams. This is shown by the reload curve
(point 4 to point 5) in figure 12, which has two phases: (i) the
initial steep portion of the curve that relates to the stiffness of
the beam with a closed crack; and (ii) the subsequent non-
linear section of the curve that reflects a gradual reduction in
stiffness due to the opening and further extension of the crack.
The start of phase two, denoted the ‘transition load’ in
figure 12, may be associated with the point at which the
bending stress in the bottom fibre of the beam overcomes
the corresponding compressive stress from the tendon i.e. the
point at which the crack starts to re-open.
At point 5, the stress applied to the polymer SP can again
be calculated from equation (1), using M, xc and Z values
appropriate to point 5. With an applied load of 4.27 kN, SP is
equal to 52.1 Nmm−2. This is still well within the elastic zone
Figure 13. Stress block diagrams for (a) points 1, 2 and 5 on figure4.35; (b) activation of the polymer tendon (point 4 on figure 4.35).T=tension, C=compression, SP=Stress applied to polymertendon, Xu=depth of neutral axis (un-cracked section),Xc=depth of neutral axis (cracked section), FP=shrinkage forceapplied by the polymer.
8
Smart Mater. Struct. 27 (2018) 075016 O Teall et al
of the polymer, so it can be assumed that yielding has not
occurred.
Upon unloading (point 6), the CMOD did not return to its
value after activation (point 4), which can again be attributed
to increased misalignment between crack faces during the
second phase of cracking.
The results indicate that the compressive stress produced
by the tendons was sufficient to close the induced crack width
by up to 85%. This crack closure is also shown by the strain
images in figure 14 produced by the DIC camera system.
The restrained shrinkage stress generated by the polymer
tendons inside the beams was calculated from the reload
curves (point 4 to point 5 on figure 12) of the PET beams after
activation.
The transition load that occurs between load points 4 and
5 is used to compute an estimate of the activated tendon stress
by assuming that the extension of the tendon due to beam
flexure is negligible at this point and that the compressive
stress in the bottom fibre of the beam caused by the pre-stress
exactly balances the bending stress from the applied moment.
This condition is expressed mathematically in equation (2).
s s=
⋅+
⋅ ⋅ ⋅+
⋅( )
A
A
A e y
I
M y
I0 2
p p p p
u u
t
u
in which Mt is the mid-span moment associated with the
applied ‘transition load’ Pt and subscript u denotes un-
cracked properties
Figure 14. Digital image correlation (DIC) camera snapshots showing closure of concrete crack using PET tendons. Axis indicates strain.(a) Loaded to 0.5 mm CMOD; (b) unloaded; (c) post-activation.
Table 6. Crack closure of unreinforced beams by PET tendon activation.
Sample
CMOD
before (mm)
CMOD
after (mm)
CMOD
closure (%)
Microscope
before (mm)
Microscope
after (mm)
Microscope
closure (%)
PET A 0.231 0.035 85 0.169 0.028 83
PET B 0.195 0.045 77 0.132 0.045 66
PET C 0.2 0.060 70 0.165 a
a
Measurement not taken due to equipment error.
Table 7. Restrained shrinkage stress of polymer tendons withinbeams—calculated values.
Sample
Polymer compressive
load (kN)
Calculated polymer
restrained shrinkage
stress (MPa) CoV (%)
PET A 1.7 19.71
PET B 1.3 15.1 11.5
PET C 1.4 16.23
Figure 15. Load-CMOD curves for PET A-r and CONT B-r samples.
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Smart Mater. Struct. 27 (2018) 075016 O Teall et al
Table 7 shows the calculated restrained shrinkage stress
values for the three PET samples using this method.
The calculated values are similar to the final stress values
observed in the tendon experiments undertaken outside of
concrete beams, given in table 4. PET B and C exhibited
lower shrinkage stress results than PET A, which agrees with
the lower percentage of crack closure observed in table 6. The
coefficient of variance of 11.5% within the beam samples
compared to 3.6% outside indicates increased variability of
results when activating the tendons inside concrete. The
concrete beam arrangements add several new variables,
including the eccentricity of the tendons and any expansion
and contraction of the concrete, which affect the final
shrinkage stress observed from the tendons.
5.4. Reinforced beams
Figure 15 shows the load-CMOD curve for the PET A-r test
beam compared to CONT B-r. This control sample was used
for comparison as it was left for an hour following loading to
investigate the effect of the steel reinforcement in closing the
crack over time.
The inclusion of the un-activated PET tendon had no
noticeable impact on the load at which concrete cracking took
place (point 1). However, there is an apparent increase in
overall stiffness of the PET beam prior to activation as shown
by a reduced CMOD reading at the maximum load of 15 kN
(point 2). On average, the CMOD of the PET beams at 15 kN
before activation was 10.6% lower than the control beams
(0.202 mm compared to 0.226 mm).
Upon activation (point 3 to point 4), the CMOD reduced
by between 26% and 39% for the PET beams, compared to an
11% reduction due to the action of the steel in CONT B-r
when left for an hour after loading. The microscope mea-
surements show very similar results, with 25%–39% closure
in PET A-r and PET B-r samples compared to 13% in CONT
B-r. Table 8 shows the crack closure upon activation for all
beams based on the CMOD and microscope measurements. A
notable exception is PET C-r which shows only 10% closure
based on microscope measurements compared to 26% based
on CMOD. The reason for this discrepancy is unclear,
although the residual crack width measured from microscope
images in this sample were the smallest of all samples,
making changes in crack width more difficult to measure.
The reload curve following activation of PET A-r (point
4 to point 5 in figure 15) again shows an apparent increase in
stiffness, as the CMOD at 15 kN load reduced from 0.194 mm
to 0.190 mm. In contrast, the CMOD of CONT B-r at 15 kN
was 0.243 mm upon reloading, an increase from 0.231 mm on
initial loading.
There appeared to be some interference with the elec-
trical activation system by the steel reinforcement. It is sug-
gested that this may be due to induced current in the steel
running parallel to the heating coils. The target internal
thermocouple temperature of 90 °C was not achieved in any
of the reinforced samples, even when the temperature on the
heating wire was raised to 150 °C. The wire temperature was
not increased beyond this point to avoid locally melting the
Table 8. Crack displacements and closure from CMOD and microscope measurements (in mm)—microscope measurements are an averageover both sides of the beam samples.
PET A-r PET B-r PET C-r CONT A-r CONT B-r CONT C-r
CMOD Micro. CMOD Micro. CMOD Micro. CMOD Micro. CMOD Micro. CMOD Micro.
Peak load 0.194 0.120 0.217 0.126 0.196 0.120 0.246 0.168 0.231 0.132 0.201 0.126
Unloaded 0.050 0.023 0.064 0.028 0.063 0.021 0.080 0.043 0.077 0.032 0.059 0.024
Activated 0.030 0.014 0.048 0.021 0.047 0.019 — — 0.069a 0.028 — —
Crack closure 0.019 0.009 0.017 0.007 0.016 0.002 — — 0.008 0.004 — —
Crack clo-
sure (%)
39% 39% 26% 25% 26% 10% — — 11% 13% — —
a
CONT B-r left for an hour following unloading, closure due to elastic effect of steel reinforcement.
Table 9. Maximum recorded temperatures during activation.
PET A-r PET B-r PET C-r
Thermocouple External Internal Wire External Internal Wire External Internal Wire
Temperature (°C) 82.7 84.6 135.8 60.8 74.2 147.0 59.2 68.1 150.0
Figure 16. Load-CMOD reload curves for PET A-r and CONT B-r.
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Smart Mater. Struct. 27 (2018) 075016 O Teall et al
PET filament. The maximum recorded temperatures are
shown in table 9.
A comparison of three successive reload cycles of PET
A-r and CONT B-r is shown in figure 16. The CMOD and
microscope measurements at peak load and after unloading
for all samples are displayed in table 10. While all samples
experienced increasing displacement at peak load with repe-
ated loading, this effect was reduced in the PET samples. On
average, the peak load displacement increased by 4.8% over
three reload cycles for the PET beams, compared to 6.6% for
CONT B-r. Moreover, the average final, unloaded CMOD
reading after all load cycles for the PET beams was
0.045 mm, 39% smaller than the 0.074 mm measured on
CONT B-r. Considering only PET A-r, which experienced the
most successful polymer activation, this reduction in final
CMOD reading increases to 54%. This suggests that the
compressive stress produced by the polymer tendon can
reduce gradual displacement over time due to cyclic loading
and that the compressive force generated by the PET was not
lost during reload cycles.
These results are similar to those produced by Yachuan
and Jinping (2008) using SMA wires in their super-elastic
form in concrete beams, in which deflections were recovered
over repeated load cycles. However, the SMA in this form
can only recover displacements, not produce a compressive
stress on the crack faces required to enhance autogenous
healing. An autonomic healing response was instead intro-
duced using brittle, adhesive-filled fibres.
Figure 17 shows DIC images of the test on sample PET
A-r. Three separate cracks are created by the loading process
in both the PET and control beams due to the presence of the
reinforcement. The cracks are still present in the sample fol-
lowing activation, and the crack closure is not as clear from
the DIC strain images as in the unreinforced samples.
Due to the steel reinforcement, the induced cracks are
much smaller in the reinforced PET samples upon unloading
(average of 0.059 mm CMOD compared to 0.209 mm CMOD
in unreinforced samples). In addition to the challenges of
achieving 90 °C across all of the filaments during activation,
the polymer tendon is also working against the misalignment
Table 10. CMOD and microscope crack measurements on reload cycles for all samples (mm unless stated).
PET A-r PET B-r PET C-r CONT A-r CONT B-r CONT C-r
CMOD Micro. CMOD Micro. CMOD Micro. CMOD Micro. CMOD Micro. CMOD Micro.
R1 peak 0.190 0.108 0.223 0.118 0.204 0.114 0.263 0.173 0.243 0.168 0.215 0.140
R1 unload 0.032 0.012 0.052 0.024 0.050 0.018 0.079 0.047 0.074 0.038 0.061 0.035
R2 peak 0.196 0.229 0.213 0.274 0.252
R2 unload 0.033 0.052 0.050 0.080 0.074
R3 peak 0.199 0.232 0.216 0.259
R3 unload 0.034 0.052 0.050 0.074
Increase over 3
cycles
4.7% 4.0% 5.9% 6.6%
Figure 17. Digital image correlation (DIC) snapshots (a) activationof tendon in PET A-r (b) loading of CONT A-r. Axis indicates strain.
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Smart Mater. Struct. 27 (2018) 075016 O Teall et al
of the crack faces and any debris wedged in the crack fol-
lowing initial loading to achieve any crack closure. In the
unreinforced samples, the residual cracks from initial loading
were large enough to allow some crack closure to occur
before effects of misalignment or debris could become
significant.
Given the limited success of these tendons within rein-
forced beams, it is proposed by the authors that this SMP
system may be better employed as full or partial replacement
for steel reinforcement, with the tendons acting both as
reinforcement and as a partial delayed prestressing system for
crack width reduction. In this way, the SMP could produce a
low level pre-stress on the concrete member to restrict crack
widths to below the levels required by current design stan-
dards, while enhancing the autogenous healing where water is
present. This approach could also reduce material costs by
replacing the expensive steel bars and avoid issues with the
polymer activation system interacting with the steel. It is also
noted that on-going work on developing higher performance
tendons and resolving some of the present activation pro-
blems should widen the range of applicability of the system in
the future.
6. Conclusions
A shape memory PET filament tendon has been developed
which can be activated via an electrical wire system to close
cracks in concrete beam samples. Restrained shrinkage stress
experiments performed on the polymer tendons alone pro-
duced stresses of 17.79–19.19MPa upon activation and
cooling.
Tendons embedded within unreinforced and reinforced
concrete beams were cracked in three-point bending before
activating the polymer using the electrical wire system. The
stress produced in unreinforced beams resulted in crack clo-
sure of up to 85% and a significant increase in beam stiffness
on reloading. In reinforced beams, despite challenges in
achieving full activation of the polymer tendons, the mea-
sured crack closure was 26%–39% based on CMOD mea-
surements across the largest crack. Incremental increases in
CMOD during repeated loading cycles on the reinforced
beams were also reduced by the polymers.
An electrically activated SMP tendon crack closure sys-
tem for concrete, as described within this paper, has not
previously been demonstrated. Based on the results of these
experiments, it is suggested that a potential use for the SMP
system may be as full or partial replacement for steel rein-
forcement in concrete structures.
Acknowledgments
Thanks must go to the EPSRC for their funding of the
Materials for Life (M4L) project (EP/K026631/1) and to
Costain Group PLC for their industrial sponsorship of the
project and author.
ORCID iDs
Oliver Teall https://orcid.org/0000-0002-1840-3495
References
Behl M and Lendlein A 2007 Shape-memory polymers Mater.
Today 10 20–8BSI 2013 BS EN 206:2013 Concrete—Specification, Performance,
Production and Conformity (London: British StandardsInstitute)
Coates P D and Ward I M 1981 Die drawing: solid phase drawing ofpolymers through a converging die Polym. Eng. Sci. 21
612–8De Rooij M R, van Tittelboom K, De Belie N and Schlangen E 2013
Self-healing phenomena in cement based materials State-of-the-Art Report of RILEM TC 221-SHC
Government H 2013 Construction 2025 Construction 2025
Industrial Strategy: government and industry in partnership.https://gov.uk/government/publications/construction-2025-strategy
Heide N T and Schlangen E 2007 Self-healing of early age cracks inconcrete 1st Int. Conf. on Self Healing Materials (Dordrecht:Springer)
Homma D, Mihashi H and Nishiwaki T 2009 Self-healing capabilityof fibre reinforced cementitious composites J. Adv. Concr.Technol. 7 217–28
Isaacs B 2011 Self-healing cementitious materials Doctor ofPhilosophy Cardiff University
Isaacs B, Lark R, Jefferson T, Davies R and Dunn S 2013 Crackhealing of cementitious materials using shrinkable polymertendons Struct. Concr. 14 138–47
Jefferson A, Joseph C, Lark R, Isaacs B, Dunn S and Weager B 2010A new system for crack closure of cementitious materials usingshrinkable polymers Cement Concr. Res. 40 795–801
Li V C, Lim Y M and Chan Y-W 1998 Feasibility study of a passivesmart self-healing cementitious composite Composites B 29
819–27Meng Q and Hu J 2009 A review of shape memory polymer
composites and blends Composites A 40 1661–72Neville A M 2012 Properties of Concrete (Upper Saddle River, NJ:
Prentice-Hall)Schneider C A, Rasband W S and Eliceiri K W 2012 NIH Image to
ImageJ: 25 years of image analysis Nat. Meth. 9 671–5Teall O, Pilegis M, Sweeney J, Gough T, Thompson G, Jefferson A,
Lark R and Gardner D 2017 Development of high shrinkagepolyethylene terephthalate (PET) shape memory polymertendons for concrete crack closure Smart Mater. Struct. 26 4
Van Breugel K 2007 Is There a Market for Self-Healing Cement-Based Materials Proc. 1st Int. Conf. Self-healing Materials
(Noordwijk aan Zee, The Netherlands, 18-20 April 2007)
Van Den Heede P, Maes M and De Belie N 2014 Influence of activecrack width control on the chloride penetration resistance andglobal warming potential of slabs made with fly ash + silicafume concrete Constr. Build. Mater. 67 74–80
Van Mier J 1996 Fracture Processes of Concrete (Boca Raton, FL:CRC Press)
Xie T 2011 Recent advances in polymer shape memory Polymer 52
4985–5000Yachuan K and Jinping O 2008 Self-repairing performance of
concrete beams strengthened using superelastic SMA wires incombination with adhesives released from hollow fibers SmartMater. Struct. 17 025020
Yang Y, Lepech M D, Yang E-H and Li V C 2009 Autogenoushealing of engineered cementitious composites under wet–drycycles Cement Concr. Res. 39 382–90
12
Smart Mater. Struct. 27 (2018) 075016 O Teall et al